a power factor oriented railway power flow controller for...

0278-0046 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.

This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication. Citation information: DOI 10.1109/TIE.2016.2615265, IEEETransactions on Industrial Electronics

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS

Abstract—Focusing on the freight-train dominant electrical railway power system (ERPS) mixed with ac-dc and ac-dc-ac locomotives (its power factorϵ[0.70,0.84]), this paper proposes a power factor oriented railway power flow controller (RPFC) for the power quality improvement of ERPS. The comprehensive relationship of the primary power factor, converter capacity, and the two phase load currents are built in this paper. Besides, as the main contribution of this paper, the optimal compensating strategy suited the random fluctuated two phase loads is analyzed and designed based on a real traction substation, for the purposes of satisfying the power quality standard, enhancing RPFC’s control flexibility, and decreasing converter’s capacity. Finally, both the simulation and the experiment are used to validate the proposed conceive.

Index Terms—Power factor; negative sequence; power

quality; power flow controller; electrical railway power system; converter

I. INTRODUCTION

ONSIDERING the cost-efficiency, the electrical trains are

fed by the single phase grid, which are supplied from the

three phase to two phase traction transformer in electrical

railway power system (ERPS). Due to the random unbalanced

two phase loads, amount of negative sequence currents (NSCs)

along with the feeder voltage fluctuation in violent are occurred

in the utilities and ERPS itself [1], [2]. Besides, though some

Manuscript received May 5, 2016; revised July 3, 2016 and August

24, 2016; accepted September 2, 2016. This work was supported in part by the National Nature Science Foundation of China under Grant 51477046 and 51377001, and in part by the International Science and Technology Cooperation Program of China under Grant 2015DFR7 0850.

S. Hu, B. Xie, Y. Li, Z. Zhang, L. Luo, and Y. Cao, are with the College of Electrical and Information Engineering, Hunan University, Changsha 410082, China (e-mail: [email protected]; [email protected]; [email protected]; [email protected]; [email protected]; yjcao@hnu. edu.cn). (Corresponding author: Yong Li.)

X. Gao is with the Dongguan Power Supply Bureau, Guangdong Power Grid Company Ltd., Dongguan 523000, China.

O. Krause is with the School of Information Technology and Electrical Engineering, The University of Queensland, Brisbane, QLD 4072, Australia (e-mail: [email protected]).

new generation trains with PWM-based front end rectifier are

launched in Chinese railway’s rapid developing period, due to

the historic reason, many old-fashion phase controlled ac-dc

locomotives still act as the main role (occupied almost 85% of

the total railroad mileage [3]), and this status cannot be

changed in a short term. Hence, excepting NSC, reactive power

or harmonics (including low and high-order components) are

also injected into the high-voltage grid [4]; it is particularly

serious in the freight-transportation dominant ERPS mixed

with ac-dc and ac-dc-ac trains, where the PFϵ[0.70,0.84] [5].

The above issues not only imperil grid reliability and security,

but also deteriorate the power quality (PQ) of the surrounding

customers. It arouses widespread attentions of related industrial

sectors and engineers in the worldwide [6]-[8].

As the popular PQ improvement rig, static var compensator

(SVC) [9], [10], static synchronous compensator (STATCOM)

[11]-[15], active filter [16]-[21], transformer integrated power

conditioner [22]-[24], railway power flow controller (RPFC)

[5], [25]-[30], and the well-designed train-mounted front end

rectifier [31]-[33] are commonly used in ERPS. Considering

the comprehensive performance, RPFC is concerned greatly by

related departments due to its compatibility – it can, unlike the

above rigs, integrate in the secondary side of almost all kinds of

traction transformer. By rebalancing the two phase active

power, and compensating the reactive power or harmonics in

each phase independently, RPFC can deal with almost all the

main PQ problems of ERPS. Additionally, the feeder voltage’s

stability and the capacity utilization ratio of the main

transformer can also be enhanced significantly [26] , [30],

which are attractive for improving ERPS’s transport capacity

and cost-efficiency.

However, the high capacity or initial investment slowdown

RPFC’s industrial application speed. Up to now, few researches

have focused on the capacity controlling of RPFC. Benefit from

the well-designed LC branches, a novel LC coupled RPFC

(LC-RPFC) proposed in [5] can effectively reduce the

VA-capacity of its active part, because the dc-link voltage can

be reduced about 30%-40% than the conventional RPFC.

However, for resent research, the compensating strategy has to

be restricted on the “full compensation model (FCM)” in the

designing process of the LC-branches, i.e., after compensation,

A Power Factor Oriented Railway Power Flow Controller for

Power Quality Improvement in Electrical Railway Power System

Sijia Hu, Member, IEEE, Bin Xie, Yong Li, Senior Member, IEEE, Xiang Gao, Zhiwen Zhang, Longfu Luo, Member, IEEE, Olav Krause, Member, IEEE, Yijia Cao, Senior Member, IEEE

C




the primary power factor (PF) equals to 1, and the primary NSC

tends to 0, which means LC-RPFC has to bear the largest

compensating current [27]. On the other hand, the Chinese

national standard [34] indicates that, the consumer can avoid of

penalty when the primary PF≥0.9 and the 95% probability

value of the primary voltage unbalance ratio Vunb%≤2% [note:

Vunb%=V-/V+×100%; V- or V+: fundamental negative or positive

sequence voltage (NSV or PSV)]. So, a large amount of

capacity is still wasted in the FCM-designed LC-RPFC.

Besides, the achievements obtained from LC-RPFC (or RPFC

[28]) are only based on the no-popular single phase ERPS

(occupied less than 1% of the total railroad mileage in China)

though it has some advantages; while, few researches focus on

the common used two phase system, the design of the LC

branches in two phase system are much complex than that in the

single one is the main obstacle. All the above unsatisfactory

aspects should be further improved in the future study.

For maximizing blocking NSC, a new compensating strategy

was proposed in [30]. It focuses on the topic of minimizing

NSC for a given RPFC’s capacity, that is to say, it has no help

on the capacity determination in the designing stage of RPFC.

Besides, considering the short circuit capacity Sd of a traction

substation is always designed within 500-1500MVA, we found

in the practical engineering project that, after a small amount of

compensation, the standard of Vunb% can be easily achieved

than the requirement of PF, especially for V/v transformer (note:

Vunb%=1.732VNI-/Sd; VN: primary normal line voltage, I-: NSC).

That is to say, the reactive power should be confirmed to be the

main compensating target of RPFC in the ac-dc locomotive

dominant ERPS with mixed trains, the regulation of NSC, then,

degrades into the subordinate one, but cannot be neglected.

To further improve RPFC’s capacity utilization capability

and control flexibility in both designing and operating stages in

freight-train dominant ERPS, in this paper, we will focus on the

solution of the following aspects:

1) Establishing the relationship between the primary PF with

RPFC’s compensating capacity; the converter’s capacity

can be flexibly designed by adjusting the primary PF.

2) In the premise of minimizing RPFC’s capacity for a given

PF, conceiving an optimal control strategy to decreasing

NSC and NSV in a satisfactory level.

3) The proposed control strategy should not only be applied

in the simple single phase ERPS, but also in the important

common used two phase system (see Fig.1).

This paper is organized as follows, the mathematical model

of the RPFC integrated two phase ERPS is presented in Section

II. In the premise of mitigation NSC, as the main contribution

of this paper, Section III gives the PF oriented optimal

compensation strategy for RPFC. Simulation and experiment

are given in Section IV and V. Section VI is the conclusion.

II. GENERAL MATHEMATICAL MODEL OF RPFC

INTEGRATED IN TWO PHASE ERPS

First, we define the frame-ABC by the V/v transformer’s

primary three phase voltage VA, VB, and VC, i.e., Frame-ABC:

A p B p C p0 , 120 , 240V V V V V V (1)

where Vp is the root mean square (RMS) value of VA, VB, and

VC.

Reference to Fig. 1, the phasor diagram of the V/v

transformer based ERPS can be obtained, as shown in Fig. 2.

From Fig. 2, we define the PF in phase-A, B, and C, i.e.,

PFA~PFC as:

A a B b C cPF cos ,PF cos ,PF cos (2)

where, φk>0 means that the current lags the voltage, otherwise,

the current leads the voltage (k=a, b, c).

It can be seen from Figs. 1 and 2 that, the output currents Iα

and Iβ of the V/v transformer in frame-pαqα and frame-pβqβ (see

Fig.2) can be expressed as

( ) ( )

( ) ( )

p q

p q

L c L p c p L q c q

I I

L c L p c p L q c q

I I

I I j I I

I I j I I

I I I

I I I, (3)

where subscript “p” and “q” represents the active and reactive

component of the corresponding variable in frame-pαqα or

frame-pβqβ, respectively.

Additionally, Fig. 2 also shows that the relationship of the p,

q components of Iα and Iβ in frame-pαqα and pβqβ satisfy:

Ai Ci Bi

i i

iLi Li ci ci v

v

Phase -α Phase -β

RPFC IsolationTransformer

Phase - APhase - BPhase - C

110kV or 220kV High - voltage Grid

V / v Transformer

27.5kV 27.5kV

dcv

kL kL

:1N :1N

Fig. 1. The typical RPFC integrated two phase ERPS (V/v transformer is adopted as the main transformer).

La

L

q

p

AV

o0

BV

CV

I

I

I

c

a

bV

V

p

q

cI

LI

LI

cI

c pI

c qI

qI

pI

c pI

c qI

qI

pI

Fig. 2. The phasor diagram of the V/v transformer based ERPC with RPFC.




tan ( ) tan

tan ( ) tan

q p L p c p

q p L p c p

I I I I

I I I I

, (4)

where a

b 120

.

Note: for V/v transformer, ∆α=30°, ∆β=90° [27].

Ignoring the converter’s losses, and assuming Vα=Vβ, the

active power balance of the back-to-back converter can lead the

result of:

c p c pI I . (5)

On the other hand, Fig. 2 indicates that Iγ’s phase angle Θγ in

frame-ABC satisfy

c c120 or tan =tan(120 ) . (6)

Based on the Kirchhoff's law, Iα, Iβ, and Iγ in frame-ABC

IαABC

, IβABC

, and IγABC

satisfy ABC ABC ABC

I I I I , (7)

where ABC

ABC

I I

I I. (8)

Substituting (3), (4), and (8) into (7), the real and imaginary

part of -Iγ, Term-I and Term-II, can be calculated as

Term-I cos sin cos sin

Term-II sin cos sin cos

p q p q

p q p q

I I I I

I I I I

(9)

Substituting (9) into (6), and considering the expressions of

Iαp, Iαq, Iβp, and Iβq in (3)-(5), the relationship of Icαp with the two

phase load active currents ILαp and ILβp can be calculated as

1 2

1 2 1 2

c p L p L p

x xI I I

x x x x

, (10)

where

1

2

sin cos tan

cos tan sin

x

x

,

c

c

120

120

.

Re-substituting (10) into (3)-(5), the compensating currents

of RPFC can be obtained as

[tan (1 ) tan ] tan

tan [tan (1 ) tan ]

c p L p L p

c p L p L p

c q L L p L p

c q L p L L p

I I I

I I I

I I I

I I I

. (11)

Multiplying the feeder voltage Vα or Vβ in the two sides of

(11), RPFC’s compensating power in phase α and β, i.e., Pcα,

Qcα and Pcβ, Qcβ, can be calculated as

[tan (1 ) tan ] tan

tan [tan (1 ) tan ]

c L L

c L L

c L L L

c L L L

P P P

P P P

Q P P

Q P P

(12)

where PLα and PLβ are the load’s active power in phase α and β.

It can be seen from (12), because ∆α, ∆β can be pre-obtained

for a certain type of a transformer (e.g., the V/v transformer and

other kind of the balance transformers [35], [36]), μα and μβ are

only determined by PFA~PFC or φa~φc [see (10) and (2)]. Hence,

the active and reactive power of the RPFC can be flexibly

adjusted by controlling the primary three phase power factors,

if the PFs of the two phase loads are pre-calculated [see φLα and

φLβ in (12)], which will be discussed later on.

III. COMPENSATING STRATEGY DESIGN

A. The Possible Compensating Scheme

For the consideration of designing convenience and the

requirement of PF≥0.9, we let

a b c

*

kPF cos [0.9,1], k=a, b, c

, (13)

where PF* is the primary reference power factor.

It can be observed from Fig. 2 that Iα, Iβ, and Iγ (or IA, IB, and

IC) may leads or lags VA, VB, and VC, respectively, which

means eight (i.e., 8=23) possible combination models with

positive or negative value are existed in φa, φb, and φc. Besides,

Fig. 2 also indicates the reactive power of converter-α is larger

than the one generated by converter-β (i.e., Icαq>Icβq), to reduce

the VA-capacity of converter-α, Iα has to be restricted lagging

than VA (i.e., φa>0), so the above eight possible combination

models of φa~φc will degenerate into four valuable candidates,

which are listed in Table I (i.e., Model-2 to -5).

B. Compensating Capacity Analysis

The VA-capacity SRPFC of the RPFC is:

converter converter

2 2 2 2

RPFC c c c c

S S

S P Q P Q

(14)

Substituting (12) into (14), the RPFC’s VA-capacity in the

five compensating model listed in Table I are shown in Fig. 3

[PLα and PLβ are the two phase loads’ active power, PF*=0.95,

and the two phase loads’ PF=0.8 (from a substation’s data)].

It can be seen from Fig. 3(a) that, the VA-capacity of RPFC

belongs to five different surfaces in Model-1~5 respectively.

The maximum SRPFC occurs in the single phase loaded

condition, in which Model-1, 2, and 4 correspond to PLα≠0,

PLαβ=0, while the opposite situation belongs to Model-3 and 5.

Additionally, a surface spliced by the surfaces of Model-2, 4,

and 5 has the minimum SRPFC. Compared with Model-1, i.e.,

FCM, the capacity decreasing ratio of this spliced surface is

about 30%, which can make the converter have higher system

reliability and efficiency. So, it can be selected as the optimal

compensating surface. If PF*=0.95, from Fig. 3(b) the optimal

TABLE I COMPENSATING SCHEME OF RPFC

*

Compensating model φa φb φc

Model-1(i.e., FCM) 0 0 0

Model-2 >0 <0 >0

Model-3 >0 <0 <0

Model-4 >0 >0 >0

Model-5 >0 >0 <0

* φk>0 (or <0) means the inductive (or capacitive) PF (k=a, b, and c).




compensating strategy (OCS) can be preliminary expressed as

* 0.95

0

0.

Model-5, MW 0.55

| : Model-4, 1.67

Model-2, 1.67 8MW

55

L L

L L L

L L

P P

P P P

P P

PFOCS . (15)

C. The NSC Mitigation Ability Analysis

Excepting of compensating reactive power, mitigation of the

NSC is another purpose of RPFC. That is to say, a satisfactory

compensating strategy should not only minimize SRPFC, but also

has the responsibility to reduce NSC within a satisfactory level.

Combing (7)-(8), the primary positive and negative sequence

currents, I+ and I-, can be deduced by ABC

2

ABC

2

ABC

11

3 1

( )(1 tan ) ( 30 )

( 30 )3

( )(1 tan ) ( 90 )

( 90 )3

L p L p

L p L p

N

I I j

N

I I j

N

II

II

I

(16)

where ξ=∠120°, N=VsN/VfN is the turn’s ratio of the main

transformer (VsN and VfN are the grid and feeder normal line

voltage respectively; as shown in Fig. 1).

From (16), the current unbalance ratio Iunb (Iunb=I-/I+) can be

obtained as follows. 2 2 2 2

a b

unb 2 2 2 2

a b

cos cos 2 cos cos cos( 180 )I

cos cos 2 cos cos cos( 60 )

(17)

From (13), (17), and Table I, the relationship of Iunb and PF*

of Model-1~ 5 are shown in Fig. 4. From Fig. 4 and Fig. 3 (a),

we can observe that, though the capacity surfaces of Model-3

and 5 are very close [Fig. 3 (a)], the NSC suppressing ability of

Model-3 is better than that of Model-5 (Fig. 4). It indicates that,

if Model-5 is substituted by Model-3, RPFC can get the better

NSC suppressing ability with almost has the same VA-capacity

of Model-5. That is to say, the compensating strategy combined

of Model-2, 4, and 3 has higher comprehensive performance

than the one combined by Model-2, 4, and 5. So the genuine

OCS should be modified from Fig. 3(b) into Fig. 5, and its

specification is given in (18).

* 0.95

0

0

Model-3, MW 0.415

| : Model-4, 1.67

Mode

.55

l-2, 1.67 8MW

L L

L L L

L L

P P

P P P

P P

PFOCS (18)

Fig. 6 gives the slopes of line OA and OB, i.e., KOA and KOB

in different PF* (note: OA and OB are the boundaries of the

three compensation model shown in Fig. 5; the loads’ PF are

Model-1

Model-3

SR

PF

C [

MV

A]

PLβ [MW] PLα [MW]

Model-2

Model-4Model-5

X: 8

Y: 0

Z: 13.84

YX

Z

(a)

PLβ [MW]

PL

α [

MW

]

Model-4

Model-5

Model-2

X: 8

Y: 0

Z: 10.1

X: 8

Y: 8

Z: 10.1

X: 8

Y: 4.4

Z: 9.976

X: 4.8

Y: 8

Z: 8.747

X: 0

Y: 8

Z: 8.061

PLβ

View directionPLα

X

Y

A

B

O

U

V

W

EC

D

(b) Fig. 3. The relationship of SRPFC with the two phase loads’ active power in the five valuable compensating models. (a) The surfaces of SRPFC with the two phase loads’ active power. (b) The xoy-projection of the surfaces in Fig. 3(a).

0.9 0.91 0.92 0.93 0.94 0.95 0.96 0.97 0.98 0.99 10

0.2

0.4

0.6

0.8

1

Model-3

*PF

unbI

Model-2

Model-5

Model-4 ≈Model-1

Fig. 4. The curves of Iunb v.s. PF* of Model-1~5.

View direction

Model-4

Model-3

Model-2

X: 8

Y: 0

Z: 10.16

X: 8

Y: 8

Z: 10.1

X: 8

Y: 4.2

Z: 10.12

X: 4.8

Y: 8

Z: 8.747

X: 0

Y: 8

Z: 8.061PLβ PLα

PLβ [MW]

PL

α [

MW

]

A

B

O

X

Y

U

V

W

EC

D

Fig. 5. The optimal compensating strategy of considering the NSC suppressing ability.




still confirmed to be 0.8, because the power factor fluctuates in

a small rang around 0.8 in the measured substation).

It can be observed from Fig. 6 that, KOA’s fluctuation

amplitude is 0.114, while, it varies in relatively large range for

KOB. For implementation of the proposed OCS, a satisfactory

performance can also be obtained by fixing KOA on 0.5, and

adjusting KOB by PF* according the blue curve shown in Fig. 6.

It can be pre-embedded in the digital controller’s memory space

in practical application.

D. Negative Sequence Standard Consideration

From (16), the primary NSC I- can be calculated as:

2 2

1 2

1( )

3L p L pI I I

N , (19)

where

1

2

[tan cos sin ] [cos( 30 ) tan sin( 30 )]

[tan cos( 30 ) sin( 30 )] [tan sin cos ]

.

The negative sequence capacity S- in the primary side is

2 2

1 23 ( ) ( )sN L L L LS V I P P K P P . (20)

Considering the Chinese national standard of the negative

sequence component is

unb

d

V 2%V

V S

V S

, (21)

where V- and V+ are the primary negative and positive voltages,

Sd is the short circuit capacity of the traction substation.

The negative sequence requirement of the proposed system

can be calculated by combining (20) and (21), i.e.,

d( ) 2%L LK P P S . (22)

Fig. 7 gives the two phase loads’ distribution chart of a real

V/v transformer based traction substation (see Table II). The

statistic results of Fig. 7 indicate that almost 95.2% of the load

points are located in the rectangle area of CEDO, where the

probability of the points distributed in ∆ACO and ∆ABO (or

∆AB1O, or ∆AB2O) is about 85%. Furthermore, we can also

find that, exceeding 50% of the load points are located on the

line OC and OD (note: some points are overlapped on these two

lines), which means the V/v transformer’s capacity utilization

ratio can be further improved in a large potential. Based on the

above statistic results, our attention should be focused on the

loads located in CEDO and its boundaries.

The surface of S- vs. PLα and PLβ (within rectangle area of

CEDO shown in Fig. 7) can be obtained based on (20) and the

Sd given in Table II, which is shown in Fig. 8. From the shape of

the surfaces shown in Fig. 8, it can be concluded that the

maximum S- of Model-3, 2, and 4 occurs on the point A, B, and

E for any given PF*, respectively.

Fig. 9 gives the relationship of the S- in A, B, and E, i.e., the

maximum S-, S-max

, with PF* for this traction substation in

Model-2~4. Obviously, the S- blocking capability of Model-4 is

much better than that of Model-3 and 2, though the latter’s S-max

0.9 0.91 0.92 0.93 0.94 0.95 0.96 0.97 0.98 0.990

0.5

1

1.5

2

2.5X: 0.9

Y: 2.326X: 0.91

Y: 2.133 X: 0.92

Y: 1.985 X: 0.93

Y: 1.856 X: 0.94

Y: 1.758 X: 0.95

Y: 1.674X: 0.96

Y: 1.613X: 0.97

Y: 1.581

X: 0.99

Y: 1.603

X: 0.9

Y: 0.475X: 0.91

Y: 0.4591

X: 0.92

Y: 0.4749

X: 0.93

Y: 0.489

X: 0.94

Y: 0.5079

X: 0.95

Y: 0.52

X: 0.96

Y: 0.533

X: 0.97

Y: 0.5401

X: 0.98

Y: 0.5581

X: 0.99

Y: 0.5729

X: 0.98

Y: 1.578

*PF

OAK

OBK

Fig. 6. The curves of slope-AO (i.e., KOA) and BO (i.e., KOB) v.s. PF

*.

Y

PLβ [MW]

PLα

[M

W]

O

A

B

1B

2B

*PF = 0.95

*PF = 0.98

Lower boundary

* PF = 0.9Upper boundary

C

D

X

E

Fig. 7. The two phase load’s distribution of a real V/v transformer based traction substation.

TABLE II THE SPECIFICATION OF A REAL V/V TRANSFORMER

Grid line voltage 110kV

Transformer Capacity

20MVA

phase-α: 10MVA

phase-β: 10MVA

Sd of the traction substation 486MVA

Short circuit impendence phase-α and β: 10%

Turn’s ratio 110kV:27.5kV

02

46

8

0

2

4

6

80

2

4

6

8

LP [MW]

LP [MW]

S

[MV

A]

A

EB

Model-4 15S 10

Model-2S

Model-3S

Model-3S (A)

Model-4 15S (E) 10

Model-2S (B)

Fig. 8. The relationship of S- with PLα and PLβ (note: PF

*=0.95).




decreases when PF* becomes large. Fig. 9 also shows that the

maximum negative sequence powers controlled by OSC are

less than the permission value 9.72MVA (i.e., 486MVA×2%),

which means the Chinese national standard can be satisfied

when PF* is set within 0.9 to 0.99. It should be remarked here is

that, if the permission line of S- crosses with other maximum S-

line of Model-2 or 3 shown in Fig. 9, the right hand abscissa of

that intersection point should be selected as the valuable PF*,

because the left one will lead Vunb out of the limit.

The capacity utilization capability of RPFC should also be

included in our concerning scope. From Fig. 10, the maximum

SRPFC’s (i.e., SRPFCmax

) reducing ratio decreases heavily when

PF*>0.95 [note: the maximum SRPFC point in PLα-PLβ panel (i.e.,

Fig. 7) is labeled in Fig. 10]. Besides, RPFC’s designing

capacity SRPFCdesign

’s decreasing ratio also shows relatively

large value (>23.43%) when PF*ϵ[0.9,0.95], it increases when

PF*→0.9 [note: ① SRPFC

design=2×max{Sconveter-α, Sconveter-β}, this

is because IGBT is a voltage sensitive device and the dc-link

voltages of converter-α and β are the same; ② Eα in Fig.10

means the maximum converter capacity belongs to converter-α

located in point E]. Considering cost-efficiency, PF* can be

selected from 0.9 to 0.95 for this traction substation.

E. Control Strategy Realization

The control system of the RPFC is plotted in Fig. 11. Some

specifications should be made for it: The FFT method or the

instantaneous reactive power theory [37] can be used for the

calculation of the load’s active and reactive power in the “PQ

block”, while the proportional resonant regulator (PS) is

adopted as the current controller for its good tracing ability in

single phase system. For the stabilization of vdc in the back-to

-back system, instead of the calculated Pcα, the real Pcα is

generated by the dc-link voltage PI controller in converter-α.

In addition, more attention has to be paid on the realization of

the “compensating power calculation” block, and the following

four steps can help us to get the target:

1) According the measured two phase loads (e.g., Fig. 7), Sd,

and the presented slops of OA and OB shown in Fig. 6, the

PF*’s regulating range can be determined for the purposes

of satisfying the negative sequence’s standard (e.g., Fig. 9)

and having relatively small capacity (e.g., Fig. 10).

2) Based on the pre-set PF* (e.g., PF

*ϵ[0.9,0.95]), the slopes

of OA and OB can be determined from Fig. 6.

3) The compensating model of OCS can be determined by the

load point’s location in the load distribution panel shown

in Fig. 5 or 7, which can be deduced by detecting the two

phase loads’ active power PLα, PLβ, and the slops of OA and

OB pre-obtained in step 2.

4) If the compensating model is obtained from step 3, φa, φb,

and φc can be calculated from Table I and (13), so as μα and

μβ [see (10)]. Hence, the compensating active and reactive

power of RPFC can be finally obtained from (12), [note: in

(12), φLα=arctan(QLα/PLα), φLβ=arctan(QLβ/PLβ)].

IV. SIMULATION

To validate the proposed OCS, the simulation model of the

studied system shown in Fig. 1 has been established. The

parameters of the main transformer, isolation transformer (IT),

and converter are listed in Tables II and III.

Fig. 12, Table IV and Fig. 13, Table V are the simulation

results in two cases. Fig. 12 corresponds the variable PF* with

0.9 0.91 0.92 0.93 0.94 0.95 0.96 0.97 0.98 0.99

0

2

4

6

8

10

*PF

max

[]

S

MV

A ( )

Moled-2S B

( )

Moled-3S A

( )

Moled-4S E

S Permission value : 9.72MVA

Fig. 9. The relationship of the primary maximum negative capacity with PF

* in Modle-2, 3, and 4.

0.9 0.92 0.94 0.96 0.98 10

5

10

15

20

25

30

35max

RPFCS 's decreasing ratio

design

RPFCS 's decreasing ratio

E EA

E

E E

C

C

C

C

E E EE

E

E

E

E

E

E

*PF

%[]

Fig. 10. The relationship of SRPFC

max’s reducing ratio [1-maximum

SRPFC/FCM based maximum SRPFC] and SRPFCdesign

’s reducing ratio [1- SRPFC

design /FCM based SRPFC

design] with PF* under the control of OCS.

Li L

i v

LP LQ

i

PQ

Compensating Power

Calculation

*PF

cP cQ PLL

V

c pI c qI

0sin( )c mI t

t

*ci

PS

PWM

cv

PQ

Compensating Power

Calculation

*PF

cQ PLL

V

c pI c qI

0sin( )c mI t

t

*ci

PS

PWM

cv

i v

LP

LP LQ

LP

cP

2 22 2c m c p c qI I I 2 22 2c m c p c qI I I

0 arctan( / )c q c pI I 0 arctan( / )c q c pI I

Grid

Load - βLoad -α

RPFCci c

i

vv

*dcV PI

dcv

Fig. 11. The control system of the OCS based RPFC.

TABLE III THE PARAMETERS OF THE ISOLATION TRANSFORMER AND RPFC

The VA-capacity of IT 5MVA

Short circuit impendence of IT 21%

IT’s turn’s ratio a 27.5kV:27.5kV

The dc-link voltage of RPFC 51.15kV

a For discussion convenience, the turn’s ratio of IT is set to be 1:1 in the

simulation model, though it is designed to be 27.5kV/1~3kV in the industrial

system, where IT has multi-secondary windings and it acts as the interface for

the small-rating back-to-back converter unit parallel connection [25]-[26] (note: the multi-level topology is unreliable for RPFC, because of it has the

risk of short circuit between the back-to-back converter units [38]).




TABLE V ACTION SEQUENCE OF THE CASE SHOWN IN FIG. 13

Time Load condition PF* Compensation

model

0.0-0.2S PLα=0MW, QLα=0Mvar;

PLβ=8MW, QLβ=6Mvar No RPFC -

0.2-0.4s PLα=0MW, QLα=0Mvar; PLβ=8MW, QLβ=6Mvar

1 FCM 0.4-0.6s PLα=8MW, QLα=6Mvar;

PLβ=8MW, QLβ=6Mvar


0.8-1.0s PLα=0MW, QLα=0Mvar;

PLβ=8MW, QLβ=6Mvar

0.95

Model-2


Model-4

1.2-1.4s PLα=8MW, QLα=6Mvar;

PLβ=0MW, QLβ=0Mvar Model-3

0 0.2 0.4 0.6 0.8 1-150

-100

-50

0

50

100

150

Time [s]

[A]

Ai Bi Ci

(a)

0 0.2 0.4 0.6 0.8 10.85

0.9

0.95

1

Time [s]

*PF

PF

(b)

0 0.2 0.4 0.6 0.8 10

0.5

1

1.5

2

2.5

Time [s]

unbV %

unbI %

100

[%]

(c)

0 0.2 0.4 0.6 0.8 10

5

10

15

Time [s]

[MV

A]

Calculated

RPFCS

Measured

RPFCS

(d)

Fig. 12. The waveforms in the condition of variable PF* with constant

load. (a) Primary three phase currents. (b) PF* and PF. (c) Voltage’s

and current’s unbalanced ratio. (d) Capacity of RPFC.

constant load, while the opposite condition belongs to Fig. 13.

Figs. 12 and 13 show that, no matter the two phase loads change

or not, the primary PF shift along with PF* with the satisfactory

performance [Fig. 12(b) and Fig. 13(b)]. Additionally, iA, iB,

and iC tend to be the balanced three phase currents when PF*

became larger [Fig. 12(a) and Fig. 13(a)], which leads Vunb%≤2%

[Fig. 12(c) and Fig. 13(c)]. Under the governance of OCS, we

can also observe from Fig. 12(d) and Fig. 13(d) that SRPFC in all

kinds of working conditions are less than that in FCM, e.g., Fig.

12(d), 0.4-0.6s: SRPFC|PF*=0.95= 0.72SRPFC|PF*=1, Fig. 13(d), 1-1.2s:

SRPFC|PF*=0.95=0.73SRPFC|PF*=1; it is coincident with the

theoretical analysis stated in Section III.

TABLE IV ACTION SEQUENCE OF THE CASE SHOWN IN FIG. 12

Time PF* Compensation

model Load condition

0.0-0.2S No RPFC -

PLα=8MW, QLα=6Mvar; PLβ=0MW, QLβ=0Mvar

0.2-0.4s 0.90 Model-3

0.4-0.6s 0.95 Model-3

0.6-0.8s 0.97 Model-3

0.8-1.0s 1.00 FCM

0 0.2 0.4 0.6 0.8 1 1.2 1.4-200

-100

0

100

200

Time [s]

[A]

Ai Bi Ci

(a)

0 0.2 0.4 0.6 0.8 1 1.2 1.40.9

0.95

1

1.05

Time [s]

*PF

PF

(b)

0 0.2 0.4 0.6 0.8 1 1.2 1.40

0.5

1

1.5

2

2.5

Time [s]

unbV %

unbI %

100

[%]

(c)

0 0.2 0.4 0.6 0.8 1 1.2 1.40

5

10

15

20

Time [s]

[MV

A] Calculated

RPFCS

Measured

RPFCS

(d)

Fig. 13. The waveforms in the condition of variable load with constant PF

*. (a) Primary three phase currents. (b) PF

* and PF. (c) Voltage’s and

current’s unbalanced ratio. (d) Capacity of RPFC.




TABLE VI THE PARAMETERS OF THE EXPERIMENTAL RPFC

Item Parameter Remarks

Grid voltage 400V –

Tα or Tβ 5kVA, 400V:100V

–

ITα or ITβ 5kVA,

100V:100V –

Ls 3mH/15A Enhance the system’s inner impedance (the equivalent Sd=73.5kVA)

L 6mH/30A –

C a, Vdc*

5mF/400V, 185V

Cf=1.88μF(filtering the dc-link current’s high frequency noise ) b

IGBT 1200V/200

A

Produced by Infineon Technologies

AG

Start resistance Rdc

1Ω/4kW Limit IGBT’s start current; first switch off S10, and then switch on S10

Snubber

resistance Rs 10Ω/200W

In the pre-charge of C, first switch off

S9, and then switch on S9.

Discharge resistance Rd

10Ω/2kW When vdc>200V, the discharge circuit starts operation

a C is electrolytic capacitor. b Cf is non inductance polypropylene capacitor.

V. EXPERIMENT

A 2×5kW RPFC was built in laboratory to further validate

the proposed strategy. Fig. 14 gives the wiring diagram and the

real rig of the experimental system. The OCS is embedded in

the main controller (TMS320F2812 DSP), while 1# and 2#

slave controller (TMS320 F2812 DSP) are obligated for the

regulation of converter-α and -β (sample frequency: 6.4kHz).

HIOKI-3198 power quality analyzer is used here for data

acquisition. The system parameters are listed Table VI.

Fig. 15 and Table VII give the waveforms and the

specifications of the experimental results. In the single phase

working condition, as the increase of PF*, iA, iB, and iC tend to

be the balanced three phase waveforms [Fig. 15(a)-(c)], the

related Vunb and Iunb are decreased, and the RPFC’s capacity is

increased, as shown in Table VII. While the similar results are

also obtained in two phase working condition [Fig. 15(d)-(e)],

except iA, iB, and iC can easier to be made into balance [contrast

Fig. 15(b) and (e)]. It is also coincidence with the theoretical

analysis aforementioned in Figs. 8 and 9.

VI. CONCLUSION

This paper proposed a power factor oriented RPFC for the

power quality improvement in the common used two phase

freight train dominated ERPS. The mathematical model of the

T

Discharge circuit

8S

10

P

N

5mF

9S

C

dcR

sR dR

cv cv

1 CT(75 : 5) 2 CT

(75 : 5)

ITIT

6mHL 6mHL

T

A BC

a bciLi

ci

Li i

ci

v v

o

400V

1S

2S

3S

4S

5S 6

S 7S

3mHs L

1#Digitalcontroler

v

ci

v

ci

PWM PWM

Load LoadAxu

Adu

fC fC10S

Pre - charge circuit

Snubberresistance

Startresistance

dcR

1PT (4 :1)2PT (4 :1)

400V

100V

400V

100V

2#Digitalcontroler

Maincontroler

v Li v Li

cQ cQ

cP

dcv

3 CT(75 : 5) 3 CT

(75 : 5)

(a)

Converter -α

Converter -β

Maincontroller1#digital

controller

2#digitalcontroller

V / v transformer

IGBT

Discharge circuit

Pre - charge circuit

dcR

dR

sR

CsL

IT

IT

L

L

HIOKI - 3198

(b)

Fig. 14. Experimental system. (a) Wiring diagram. (b) Real rig.

20ms2A / div

Ai Bi Ci

(a)

20ms Ai Bi Ci1A / div

(b)

Bi CiAi20ms1A / div

(c)

2A / div20ms Ai Bi Ci

(d)

Bi CiAi20ms1A / div

(e)

Fig. 15. The waveforms in the condition of variable load with constant PF

*. (a) Primary three phase currents. (b) PF

* and PF. (c) Voltage’s and

current’s unbalanced ratio. (d) Capacity of RPFC. Experimental waveforms. (a) PLα=566W, QLα=424var; PLβ=0W, QLβ=0var; no RPFC. (b) PLα=566W, QLα=424var; PLβ=0W, QLβ=0var; PF

*=0.95. (c)

PLα=566W, QLα=424var; PLβ=0W, QLβ=0var; PF*=1. (d) PLα=362W,

QLα=271var; PLβ=362W, QLβ=271var; no RPFC. (e) PLα=362W,

QLα=271var; PLβ=362W, QLβ=271var; PF*=0.95.

TABLE VII THE SPECIFICATIONS OF THE EXPERIMENTAL RESULTS

Load condition

PF* Grid PF

Vunb% Iunb% SRPFC[VA] a

SRPFCcal SRPFC

mea

PLα=566W, QLα=424var;

PLβ=0W,

QLβ=0var

No RPFC 0.656 0.974 96.7 0 –

0.95 0.948 0.362 36.2 719.0 745.3

1.00 0.995 0.062 3.20 978.8 1015.5

PLα=362W,

QLα=271var;

PLβ=362W, QLβ=271var

No RPFC 0.701 0.644 49.1 0 –

0.95 0.941 0.112 5.30 456.9 473.6

a SRPFCcal: the calculated SRPFC; SRPFC

mea: the measured SRPFC.




RPFC integrated ERPS and the comprehensive design method

of the proposed control strategy are given in detail, based on a

real traction substation. The simulation and the experimental

results verify the correctness of the proposed conceives.

In the premise of satisfying the standards of the reactive

power and NSV, this paper gives an optimal control strategy for

the PQ improvement, control flexibility enhancement, and the

reduction of RPFC’s compensating and designing capacity in

two or single phase RPFC integrated ERPS. That is to say, this

control method can make the system have an attractive high

cost-efficiency in two or single phase traction load conditions.

REFERENCES

[1] S. Chen, R. Li, and H.4 Hsiang, “Traction system unbalance problem-

analysis methodologies,” IEEE Trans. Power Del., vol. 19, no. 4, pp. 1877–1883, Oct. 2004.

[2] J. Kilter, T. Sarnet, and T. Kangro, “Modelling of high-speed electrical

railway system for transmission network voltage quality analysis: Rail

Baltic case study,” in Proc. Elect. Power Qual. Supply Reliab. Conf. (PQ),

2014, pp. 323–328. [3] Z. He, H. Hu, Y. Zhang, and S. Gao, “Harmonic resonance assessment to

traction power-supply system considering train model in China

high-speed railway,” IEEE Trans. Power Del., vol. 29, no. 4, pp. 1735–1743, Aug. 2014.

[4] G. Raimondo, P. Ladoux, A. Lowinsky, H. Caron, and P. Marino,

“Reactive power compensation in railways based on AC boost choppers,” IET Electr. Syst. Transp., vol. 2, no. 4, pp. 169–177, Jua. 2012.

[5] Z. Zhang, Y. Li, L. Luo, P. Luo, Y. Cao, Y. Chen et al, “A new railway

power flow control system coupled with asymmetric double LC branches,” IEEE Trans. Power Electron., vol. 30, no. 10, pp. 5484–5498, Oct. 2015.

[6] S. Gazafrudi, A. Langerudy, E. Fuchs, and K. Al-Haddad, “Power quality

issues in railway electrification: a comprehensive perspective,” IEEE IEEE Trans. Ind. Electron., vol. 62, no. 5, pp. 3081–3090, May. 2015.

[7] T. Uzuka, “Faster than a speeding Bullet: An overview of Japanese

high-speed rail technology and electrification,” IEEE Electrification Mag., vol.1, no.1, pp. 11–20, Sep. 2013.

[8] M. Brenna, F. Foiadelli, and D. Zaninelli, “Electromagnetic model of

high speed railway lines for power quality studies,” IEEE Trans. Power Syst., vol. 25, no. 3, pp. 1301–1308, Aug. 2010.

[9] H. Wang, Y. Liu, K. Yan, Y. Fu, and C. Zhang, “Analysis of static VAr

compensators installed in different positions in electric railways,” IET Elect. Syst. Transp., vol. 5, no. 3, pp. 129–134, Jan. 2015.

[10] G. Zhu, J. Chen, and X. Liu, “Compensation for the negative-sequence

currents of electric railway based on SVC,” in Proc. ICIEA. Conf., 2008, pp. 1958–1963.

[11] K. Fujii, K. Kunomura, K. Yoshida, A. Suzuki, S. Konishi, M. Daiguji et

al., “STATCOM applying flat-packaged IGBTs connected in series,” IEEE Trans. Power Electron., vol. 20, no. 5, pp. 1125–1132, Sep. 2005.

[12] R. Grunbaum, J. Hasler, T. Larsson, and M. Meslay, “STATCOM to

enhance power quality and security of rail traction supply,” in Proc. 8th Int. Symp. Adv. Electro-Mech. Motion Syst. Electr. Drives, Lille, France,

2009, pp. 1–6.

[13] A. Bueno, J. Aller; J. Restrepo, R. Harley, T. Habetler, “Harmonic and unbalance compensation based on direct power control for electric

railway systems,” IEEE Trans. Power Electron., vol.28, no.12, pp.

5823-5831, Dec. 2013. [14] B. Gultekin, C. Gercek, T. Atalik, M. Deniz, N. Bicer, M. Ermis et al.,

“Design and implementation of a 154-kV ±50-Mvar transmission

STATCOM based on 21-level cascaded multilevel converter,” IEEE Trans. Ind. Appl., vol. 48, no. 3, pp. 1030–1045, May. 2012.

[15] P. Ladoux, G. Raimondo, H. Caron, and P. Marino, “Chopper-controlled

steinmetz circuit for voltage balancing in railway substations,” IEEE Trans. Power Electron., vol. 28, no. 12, pp. 5813–5882, Dec. 2013.

[16] S. Senini and P. Wolfs, “Hybrid active filter for harmonically unbalanced

three phase three wire railway traction loads,” IEEE Trans. Power Electron, vol. 15, no. 4, pp. 702–710, Jul. 2000.

[17] S. Rahmani, A. Hamadi, K. Al-Haddad, and A. Dessaint, “A combination of shunt hybrid power filter and thyristor controlled reactor for power

quality,” IEEE Trans. Ind. Electron., vol. 61, no. 5, pp. 2152–2164, May.

2014.

[18] H. Akagi and K. Isozaki, “A hybrid active filter for a three-phase 12-pulse

diode rectifier used as the front end of a medium-voltage motor drive,” IEEE Trans. Power Electron., vol. 27, no. 1, pp. 69–77, Jan. 2012.

[19] A. Bhattacharya, C. Chakraborty, and S. Bhattacharya, “Parallel

connected shunt hybrid active power filters operating at different switching frequencies for improved performance,” IEEE Trans. Ind.

Electron., vol. 59, no. 11, pp. 4007–4019, Nov. 2012.

[20] P. Tan, P. Loh, and D. Holmes, “Optimal impedance termination of 25-kV electrified railway systems for improved power quality,” IEEE

Trans. Power Del., vol. 20, no. 2, pp. 1703–1710, Apr. 2005.

[21] P. Tan, P. Loh, and D. Holmes, “A robust multilevel hybrid compensation system for 25-kV electrified railway applications,” IEEE Trans. Power

Electron., vol. 19, no. 4, pp. 1043–1052, Jul. 2004. [22] S. Hu, Z. Zhang, Y. Li, L. Luo, Y. Cao, and C. Rehtanz, “A new

half-bridge winding compensation based power conditioning system for

electric railway with LQRI,” IEEE Trans. Power Electron., vol. 29, no. 10, pp. 5242–5256, Oct. 2014.

[23] S. Hu, Z. Zhang, Y. Chen, G. Zhou, Y. Li, L. Luo et al., “A new integrated

hybrid power quality control system for electrical railway,” IEEE Trans. Ind. Electron., vol. 62, no. 10, pp. 6222–6232, Oct. 2015.

[24] S. Hu, Y. Li, B. Xie, M. Chen, Z. Zhang, L. Luo et al., “A y-d

multifunction balance transformer-based power quality control system for single-phase power supply system,” IEEE Trans. Ind. Appl., vol. 52, no. 2,

pp. 1270–1279, Mar. 2016.

[25] T. Uzuka, S. Ikedo, K. Ueda, Y. Mochinaga, S. Funahashi, and K. Ide, “Voltage fluctuation compensator for shinkansen,” Trans. IEE Japan, vol.

162, no. 4, pp. 25–34, May. 2008.

[26] M. Ohmi, and Y. Yoshii, “Validation of railway static power conditioner in Tohoku Shinkansen on actual operation,” in Proc. Int. Conf. Power

Electron., 2010, pp. 2160–2164.

[27] A. Luo, F. Ma, C. Wu, S. Ding, Z, Shuai, and Q. Zhong, “A dual-loop control strategy of railway static power regulator under V/V electric

traction system,” IEEE Trans. Power Electron., vol. 26, no. 7, pp. 2079–

2091, Jul. 2011. [28] N. Dai, M. Wong, K. Lao, and C. Wong, “Modelling and control of a

railway power conditioner in co-phase traction power system under

partial compensation,” IET Power Electron., vol. 7, no. 5, pp. 1044–1054, May. 2014.

[29] F. Ma, Q. Xu, Z. He, C. Tu, Z. Shuai, A. Luo et al., “A railway traction

power conditioner using modular multilevel converter and its control

strategy for high-speed railway system,” IEEE Trans. Transp. Electrif.,

vol. 2, no. 1, pp. 96–109, Mar. 2016.

[30] D. Zhang, Z. Zhang, W. Wang, and Y. Yang, “Negative sequence current optimizing control based on railway static power conditioner in V/v

traction power supply system,” IEEE Trans. Power Electron., vol. 31, no.

1, pp. 2079–2091, Jan. 2016. [31] B. Bahrani, and A. Rufer, “Optimization-based voltage support in traction

networks using active line-side converters,” IEEE Trans. Power Electron.,

vol. 28, no. 2, pp. 673–685, Feb. 2013. [32] C. Zhao, D. Dujic, A. Mester, J. Steinke, M. Weiss, S. Schmid et al.,

“Power electronic traction transformer—medium voltage prototype,”

IEEE Trans. Ind. Electron., vol. 61, no. 7, pp. 3257–3269, Jul. 2014. [33] D. Dujic, C. Zhao, A. Mester, J. Steinke, M. Weiss, S. Schmid et al.,

“Power electronic traction transformer—low voltage prototype,” IEEE

Trans. Power Electron., vol. 28, no. 12, pp. 5522–5534, Dec. 2013. [34] Quality of Electric Energy Supply Admissible Three Phase Voltage

Unbalance, National Standard GB/T 15543-2008, 2008.

[35] T. Kneschke, “Control of utility system unbalance caused by single-phase

electric traction,” IEEE Trans. Ind. Appl., vol. IA-21, no. 6, pp. 1559–

1570, Nov. 1985.

[36] F. Ciccarelli, M. Fantauzzi, D. Lauria, and R. Rizzo, “Special transformers arrangement for ac railway systems,” in Proc. Electrical

Systems for Aircraft, Railway and Ship Propulsion (ESARS), 2012, pp. 1–6.

[37] H. Akagi, E. Watanabe, and M. Aredes, Instantaneous power theory and

applications to power conditioning. New Jersey, USA: IEEE Press, 2007. [38] M. Miranbeigi, H. Iman-Eini, and M. Asoodar, “A new switching strategy

for transformer-less back-to-back cascaded H-bridge multilevel

converter,” IET Power Electron., vol. 7, no. 7, pp. 1868–1877, Jan. 2014.




Sijia Hu (S’14-M’16) was born in Hunan, China, in 1987. He received the B.Sc. and Ph.D. degrees in electrical engineering (and automat- ion) from Hunan University of Science and Tec- hnology (HNUST), Xiangtan, China, and Hunan University (HNU), Changsha, China, in 2010 and 2015, respectively. Since 2016, he is an Assistant Professor of Electrical Engineering with HNU. His research interests include power flow and power quality control of electric railway power systems, new topology converters, and stability and power

quality analyses and control of multi-converter systems.

Bin Xie was born in Hunan, China, in 1990. He received the B.Sc. degree in electrical engineering from Shaoyang University in 2013. He is currently working toward the Ph.D. degree of electrical engineering in the College of Electrical and Information Engineering, Hunan University, Changsha, China.

His research interests include power quality analysis and control of electric railway power system, new topology power flow controller.

Yong Li (S’09–M’12–SM’14) was born in Henan, China, in 1982. He received the B.Sc. and Ph.D. degrees in 2004 and 2011, respectively, from the College of Electrical and Information Engine- ering, Hunan University, Changsha, China.

Since 2009, he worked as a Research Associate at the Institute of Energy Systems, Energy Efficiency, and Energy Economics (ie

3),

TU Dortmund University, Dortmund, Germany, where he received the second Ph. D. degree in June 2012. After then, he was a Research Fellow with The University of Queensland, Brisb- ane, Australia. Since 2014, he is a Full Professor

of electrical engineering with Hunan University. His current research interests include power system stability analysis and control, ac/dc energy conversion systems and equipment, analysis and control of power quality, and HVDC and FACTS technologies.

Dr. Li is a member of the Association for Electrical, Electronic and Information Technologies (VDE) in Germany.

Xiang Gao was born in Hunan, China, in 1990. He received the B.Sc. and M.Sc. degrees in electrical engineering from Hunan Institute of Engineering and Hunan University in 2013 and 2016 respectively. He is currently working with Dongguan Power Supply Bureau, Guangdong Power Grid Company Ltd., Dongguan, China.

His research interests include power quality analysis and control of electric railway power system, power system protection.

Zhiwen Zhang was born in Hunan, China, in 1963. He received the B.Sc. and M.Sc. degrees in electrical engineering and Ph.D. degree in control theory and control engineering from Hunan University, China, in 1987, 1990, and 2006, respectively.

From 1992 to 1993 and 2006 to 2007, he was a visiting scholar at Tsinghua University, Beijing, China and a Visiting Professor at Ryerson University, Toronto, Canada, respectively. He is currently a Full Professor in the College of

Electrical and Information Engineering, Hunan University, Changsha, China. His research interests include power quality analysis and control

of electric railway power system, theory and new technology of ac/dc energy transform, theory and application of new-type electric apparatus, harmonic suppression for electric railway, power electronics application, and computer control.

Longfu Luo (M’09) was born in Hunan, China, in 1962. He received the B.Sc., M.Sc., and Ph.D. degrees from the College of Electrical and Infor- mation Engineering, Hunan University, Chang- sha, China, in 1983, 1991, and 2001, respectively.

From 2001 to 2002, he was a Senior Visiting Scholar at the University of Regina, Regina, SK, Canada. Currently, he is a Full Professor of electrical engineering at the College of Electrical and Information Engineering, Hunan University.

His current research interests include the design and optimization of modern electrical equipment, the development of new converter transformer, and the study of corresponding new HVDC theories.

Olav Krause (M’05) was born in Germany in 1978. He received the Dipl.-Ing. (ME.) and Dr.-Ing. (Ph.D.) degrees in electrical engineering from the TU Dortmund University, Dortmund, Germany, in 2005 and 2009, respectively.

Currently, he is a Lecturer of Electrical Engine- ering in the School of Information Technology and Electrical Engineering, The University of Queensland, Brisbane, Australia. His main research interests include distribution network auto-

mation, with a focus on state estimation under measurement deficiency and power system load ability determination. This is complemented by work on techniques of probabilistic and harmonic power-flow analysis.

Yijia Cao (M’98-SM’13) was born in Hunan, China, in 1969. He received the Graduation degree from Xi’an Jiaotong University, Xi’an, China, in 1988, and received the M.Sc. and Ph.D. degrees from Huazhong University of Science and Technology (HUST), Wuhan, China, in 1991 and 1994, respectively.

From September 1994 to April 2000, he worked as a Visiting Research Fellow and Rese- arch Fellow at Loughborough University, Liver- pool University, and University of the West Eng- land, U.K. From 2000 to 2001, he was a Full

Professor of HUST, and from 2001 to 2008, he was a Full Professor of Zhejiang University, China. He was appointed as a Deputy Dean of College of Electrical Engineering, Zhejiang University, in 2005. Currently, he is a Full Professor and the Vice President of Hunan University, Changsha, China. His research interests include power system stability control and the application of intelligent systems in power systems.

a power factor oriented railway power flow controller for...

Documents